ASM Metals HandBook Vol. 8 - Mechanical Testing and Evaluation

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Fig. 12 Barreling during a test when the friction coefficient is 1.00 at the specimen

loading face. Note that as the deformation increase, points A, B, and C originally on the

specimen sides, move to the loading face.

Friction on the loading face causes rollover. As shown in Fig. 12, points originally on the sides of the specimen

are ultimately located on the specimen end face. Use of a high-pressure lubricant at the loading surface of the

specimen reduces friction. One such material commonly used is 0.1 mm (0.004 in.) thick Teflon sheet. The

action of the lubricant may be enhanced if the bearing surfaces that apply the load are hard and highly polished.

The use of tungsten-carbide bearing blocks is recommended for all materials undergoing compression testing.

Other techniques have been used to reduce nonuniformity of stress and strain distributions along the gage

length (Ref 15, 16).

The contact area between the lateral faces of the specimen and the lateral support guides of the testing jig must

be well lubricated. Personnel engaged in sheet compression testing should become familiar with the literature

on the subject. A selected bibliography on this subject is given in ASTM E 9.

Testing Machine Capacity

In a compression test performed to large strains (e.g., to obtain fracture data), a large load capacity may be

required. For example, consider four medium-length cylindrical specimens suggested in ASTM E 9, where

specimens are specified with diameters that range from 12.7 to 28.4 mm (0.50 to 1.12 in.) and with length-to-

diameter ratios of 3. Using these specimen sizes, consider the testing of a material with a yield stress of 1380

MPa (200 ksi) and a compression strain-hardening exponent of 0.05. Figure 13 illustrates the load-capacity

requirements to reach a height reduction of 60% for each of the four cylinders recommended in ASTM E 9. The

maximum required load is approximately 3.5 times the load at yield. The required capacity for testing the same

specimens to failure at 60% strain in tension would be no more than 1.5 times the yield loads.

Fig. 13 Load requirements for compressing specimens of various diameters made of a

material with a yield stress of 1380 MPa (200 ksi) and a strain-hardening exponent of

0.05. Diameters: A = 28.4 mm (1.12 in.), B = 25.4 mm (1.00 in.), C = 20.3 mm (0.80 in.), D =

12.7 mm (0.50 in.). Length-to-diameter ratio (L/D) = 3

Medium-Strain-Rate Testing

Medium-strain-rate compression testing with conventional load frames is very similar to low-strain-rate

compression testing. For medium-rate testing, the load frames require the capability to generate higher

crosshead or ram velocities. An important consideration is the stiffness of the machine, as discussed in more

detail in the article “Testing Machines and Strain Sensors” in this Volume. For tests at a uniform strain rate, a

high machine stiffness is desired; techniques to increase the stiffness of a hydraulic machine are described in

Ref 17. This section describes some of the techniques used to obtain medium strain rates with conventional test

frames and additional experimental factors for measurement of load and strain at medium rates.

Grip design for compression testing at medium strain rates requires the same considerations that apply to grip

design for low strain rates. The compression specimen typically is sandwiched between two hard, polished

platens that are placed in a subpress designed to maintain parallel faces during deformation. A typical grip

assembly is shown in Fig. 14, in which a compression specimen (5.1 mm, or 0.2 in., long by 5.1 mm, or 0.2 in.,

diam) is in place and ready for testing. The ram is shown in position and is separated from the subpress by

approximately 20 mm (0.8 in.). This gap allows time (approximately 2 ms at the highest ram velocity) for the

ram to accelerate to the specified velocity. In this test, the stroke of the ram must be set accurately to ensure the

desired deformation.

Fig. 14 Subpress assembly for medium strain-rate testing with conventional load frame.

The specimen, which is 5.1 mm (0.2 in.) diam by 5.1 mm (0.2 in.) long, is sandwiched

between two highly polished platens. A quartz load washer is shown positioned above the

subpress assembly.

Measurement of Load and Displacement. As the strain rate increases, the measurement of load and

displacement becomes increasingly more difficult. The requirement for adequate frequency response in the

signal conditioners and the problems associated with load-cell ringing were discussed in the introduction to this

article. In this section, the measurement of load and displacement at medium strain rates is described in more

detail.

Measurement of Load. A typical load cell determines load by measuring displacement in an elastic member,

such as a diaphragm or cylinder. The displacements are measured with bonded strain gages; this gives the load

cell sufficient intrinsic frequency response for testing at medium strain rates. However, a problem often arises

due to ringing in the load cell. The load cell has a natural frequency of vibration determined by geometry and

physical properties, such as density and elastic modulus. Typical load cells have a natural frequency in the

range of 500 to 5000 Hz. In effect, the natural frequency of vibration sets the bandwidth of the load-measuring

system.

By this criterion alone, load cells should be sufficient for compression testing at strain rates as high as 100 s

-1

However, the transient response of the load cell in practice limits the measurement to much lower strain rates.

When a constant-strain-rate test is desired, deformation must be initiated by an impact due to the acceleration

time required by the ram. This impact can excite the natural vibrational mode of the load cell, which will

produce oscillations in the output signal that can mask the actual load measurement. Unless the impact is

dampened by some means, load measurement at strain rates greater than about 1 s

-1

can be subject to load-cell

ringing.

Ringing of the load cell can be minimized by selecting a load cell with a high vibrational frequency. If the

natural frequency is sufficiently high, the vibrational mode may not be excited by the impact, or if excited, it

may be possible to remove it from the signal with a low pass filter. Another method to reduce ringing is to

dampen the impact that initiates deformation within the specimen. Often, a thin layer of deformable material

placed between the impacting surfaces is sufficient to remove the higher frequencies generated by the impact

that can excite the natural frequency of the load cell. For example, in the configuration shown in Fig. 14, a

single loop, approximately 50 mm (2 in.) in diameter, of 0.51 mm (0.02 in.) diameter lead-tin solder wire

placed on the impacting face of the hydraulic ram was found to be effective in minimizing load-cell ringing.

Such layers, however, may complicate measurement of displacement within the specimen.

At strain rates close to 100 s

-1

, the standard load cell either may not possess the necessary frequency response,

or it may ring excessively. These characteristics can make the load cell inadequate for load measurement.

Under these conditions, a quartz piezoelectric device, such as a load washer (Fig. 14), is useful. The load

washer is convenient because it is easily adapted to a compression test; it also has excellent intrinsic frequency

response and a high fundamental vibrational frequency. However, these devices require special signal

conditioning and low-capacitance cables.

Measurement of Strain. The direct measurement of strain at medium strain rates presents a challenge. Many of

the devices typically used for low-strain-rate testing are inappropriate at medium strain rates. Extensometers,

for example, may have the necessary response characteristics for medium-strain-rate testing. However, it is

difficult to ensure that the rapid and large displacement in small compression specimens will not damage the

fragile extensometer.

Many hydraulic test frames use a linear variable differential transformer (LVDT) to control the motion of the

hydraulic ram. This LVDT signal is comprised of displacements within the specimen as well as elastic

displacements throughout the test frame. To relate this signal to displacements within the specimen, the latter

contribution must be subtracted; this problem also is encountered at low strain rates. If a deformable material is

placed between the impact surfaces to dampen the impact, the displacements within this layer also must be

subtracted from the LVDT signal.

A common practice is to mount the LVDT at an off-axis position adjacent to the specimen. The benefit of this

configuration is that a displacement measurement is possible between two points that are quite close to the

specimen; this measurement includes less of the elastic deformation in the load frame. When a measurement is

made at an off-axis position, it is important to verify that the measurement truly represents displacements

within the sample. Often, two LVDT units are mounted at diametrically opposite positions, and their outputs

are processed to eliminate the effects of nonplanar motion. The LVDT suffers from an intrinsic frequency-

response limitation determined by the excitation frequency. Standard excitation frequencies are in the range of

1 to 5 kHz, which limits the frequency response to around 100 to 500 Hz.

Velocity transducers, which have good intrinsic frequency response, have been used to measure the motion of

the specimen and grip assembly (Ref 17). Their output can be integrated electronically or by computer to obtain

the displacement. Generally, these also require mounting at off-axis locations.

Strain measurement by noncontact methods is becoming more common with optical extensometers or laser

interferometers. Laser interferometers, which are capable of operating at high sampling rates, can be used to

measure strain at strain rates exceeding 10

-1

Types of Compressive Fracture

For all but the most ductile materials, cylindrical specimens develop cracks when they are compressed. The

cracks generally initiate on the outer surface of the compressed specimen. As the specimen is further deformed,

the initiated cracks propagate, and new cracks form. Some different modes of compression fracture are

described in Ref 18 and some examples are described in the following sections.

Orange Peel Cracking. In many materials, roughening or wrinkling of the surface (orange peel effect) occurs

prior to compressive cracking. This effect is particularly prominent in some aluminum alloys. An extreme

example is illustrated for an aluminum alloy 7075-T6 specimen in Fig. 15. The specimen is shown after 72%

deformation. Wrinkling first appeared at 10 to 15% compressive deformation, and macrocracking occurred

after 50 to 60% deformation. Microscopic examination revealed many microcracks in the valleys of the

wrinkles, with greatest concentration in the equatorial region of the specimen. Defining a compression strength

or a strain criterion of fracture would be difficult for this material.

Fig. 15 Two views of a 72% compressed specimen of aluminum alloy 7075-T6 displaying

orange peel effect. The loading axis is vertical. Extensive macrocracking is evident in the

severely wrinkled surface. Microscopic examination of the surface revealed extensive

microcracking in the valleys of the wrinkles. Source: Ref 16

Macrocracks in Steel. A case in which macrocracks form without apparent precursor microcracks is shown in

Fig. 16. The material is AISI-SAE 4340 steel tempered at 204 °C (400 °F), yielding a hardness of 52 HRC. The

cracks initiated one at a time and extended across the surface of the specimen almost instantaneously. The first

cracks appeared when the compressive deformation reached 30%, and other cracks continued to initiate until

the test was concluded at 72% deformation, which is the condition shown in Fig. 16. The specimen was still

intact, and subsequent sectioning revealed that the cracks penetrated inward a distance of diameter.

Fig. 16 Shear cracks in a 72% compressed specimen of AISI-SAE 4340 steel. The cracks

initiated one at a time, starting when the deformation was 30%. Source: Ref 16

Microcrack to Macrocrack Coalescence. In some tungsten alloys, the first visible evidence of fracture is a shear

macrocrack that appears at the equator of the specimen after 45 to 50% compressive deformation. However,

using fluorescent-dye penetrant methods, microcrack initiation was detected at 25% deformation (Ref 18). For

this material, if crack initiation is the criterion of failure, it is necessary to state the method of crack detection

with the selected parameter for strength.

References cited in this section

13. G. Gerard, Introduction to Structural Stability Theory, McGraw-Hill, 1962, p 19–29

14. R. Papirno and G. Gerard, “Compression Testing of Sheet Materials at Elevated Temperatures,”

Elevated Compression Testing of Sheet Materials, STP 303, ASTM, 1962, p 12–31

15. T.C. Hsu, A Study of the Compression Test for Ductile Materials, Mater. Res. Stand., Vol 9 (No. 12),

Dec 1969, p 20

16. R. Chait and C.H. Curll, “Evaluating Engineering Alloys in Compression,” Recent Developments in

Mechanical Testing, STP 608, ASTM, 1976, p 3–19

17. R.H. Cooper and J.D. Campbell, Testing of Materials at Medium Rates of Strain, J. Mech. Eng. Sci.,

Vol 9, 1967, p 278

18. R. Papirno, J.F. Mescall, and A.M. Hansen, “Fracture in Axial Compression of Cylinders,”

Compression Testing of Homogeneous Materials and Composites, R. Chait and R. Papirno, Ed., STP

808, ASTM, 1983, p 40–63

Uniaxial Compression Testing

Howard A. Kuhn, Concurrent Technologies Corporation

Acknowledgments

Portions of this article were adapted from R. Papirno, Axial Compression Testing (p 55–58) and P.S.

Follansbee and P.E. Armstrong, Compression Testing by Conventional Load Frames at Medium Strain Rates (p

192–193) in Mechanical Testing, Vol 8, ASM Handbook, ASM International, 1985.

Uniaxial Compression Testing

Howard A. Kuhn, Concurrent Technologies Corporation

References

1. T. Erturk, W.L. Otto, and H.A. Kuhn, Anisotropy of Ductile Fracture—An Application of the Upset

Test, Metall. Trans., Vol 5, 1974, p 1883

2. W.A. Kawahara, Tensile and Compressive Materials Testing with Sub-Sized Specimens, Exp. Tech.,

Nov/Dec, 1990, p 27–29

3. W.A. Backofen, Deformation Processing, Addison-Wesley, Reading, MA, p 53

4. J.H. Faupel and F.E. Fisher, Engineering Design, John Wiley & Sons, 1981, p 566–592

5. G.E. Dieter, Mechanical Metallurgy, 2nd ed., McGraw-Hill, 1976, p 561–565

6. G.E. Dieter, Evaluation of Workability: Introduction, Forming and Forging, Vol 14, ASM Handbook,

ASM International, 1988, p 365

7. J.E. Hockett, The Cam Plastometer, in Mechanical Testing, Vol 8, ASM Handbook, ASM International,

1985, p 197

8. R. Herbertz and H. Wiegels, Ein Verfahren Zur Verwirklichurg des Reibungsfreien Zylinderstanch

versuchs für die Ermittlung von Fliesscurven, Stahl Eisen, Vol 101, 1981, p 89–92

9. K. Lintermanns Fander, “The Flow and Fracture of Al-High Mg-Mn Alloys at High Temperatures and

Strain Rates,” Ph.D. dissertation, University of Pittsburgh, 1984, p 258

10. H.A. Kuhn, P.W. Lee, and T. Erturk, A Fracture Criterion for Cold Forging, J. Eng. Mater. Technol.

(Trans. ASME), Vol 95, 1973, p 213–218

11. H.A. Kuhn, Workability Theory and Application in Bulk Forming Processes, Forming and Forging, Vol

14, ASM Handbook, ASM International, 1988, p 389–391

12. J.J. Jonas and M.J. Luton, Flow Softening at Elevated Temperatures, Advances in Deformation

Processing, J.J. Burke and V. Weiss, Ed., Plenum, 1978, p 238

13. G. Gerard, Introduction to Structural Stability Theory, McGraw-Hill, 1962, p 19–29

14. R. Papirno and G. Gerard, “Compression Testing of Sheet Materials at Elevated Temperatures,”

Elevated Compression Testing of Sheet Materials, STP 303, ASTM, 1962, p 12–31

15. T.C. Hsu, A Study of the Compression Test for Ductile Materials, Mater. Res. Stand., Vol 9 (No. 12),

Dec 1969, p 20

16. R. Chait and C.H. Curll, “Evaluating Engineering Alloys in Compression,” Recent Developments in

Mechanical Testing, STP 608, ASTM, 1976, p 3–19

17. R.H. Cooper and J.D. Campbell, Testing of Materials at Medium Rates of Strain, J. Mech. Eng. Sci.,

Vol 9, 1967, p 278

18. R. Papirno, J.F. Mescall, and A.M. Hansen, “Fracture in Axial Compression of Cylinders,”

Compression Testing of Homogeneous Materials and Composites, R. Chait and R. Papirno, Ed., STP

808, ASTM, 1983, p 40–63

Hot Tension and Compression Testing

Dan Zhao, Johnson Controls, Inc.; Steve Lampman, ASM International

Introduction

HIGH-TEMPERATURE MECHANICAL PROPERTIES of metals are determined by three basic methods:

• Short-term tests at elevated temperatures

• Long-term tests of creep deformation at elevated temperatures

• Short-term and long-term tests following long-term exposure to elevated temperatures

This article focuses on short-term tension and compression testing at high temperatures. The basic methods and

specimens are similar to room-temperature testing, although the specimen heating, test setup, and material

behavior at higher temperatures do introduce some additional complexities and special issues for high-

temperature testing.

Two types of long-term testing for high-temperature applications are not discussed in this article. The first type

is long-term exposure testing, where materials are exposed to high temperatures prior to mechanical testing at

either ambient or elevated temperatures. This type of testing is needed for the evaluation of metallurgical

changes that can occur during exposure to high temperatures. The second type of long-term test for many high-

temperature structural applications is the creep test. When the application temperature, T, of a stressed metallic

or ceramic material is in the range of about 0.3 T

< T < 0.6 T

(where T

is the melting point of the material

in Kelvin), the stressed material undergoes a continuous accumulation of plastic strain (i.e., creep) over time.

The continuous accumulation of creep strain can occur even when the material is stressed below its elastic

limit; therefore many high-temperature structural applications require creep testing. This type of testing is

discussed in more detail in the Section “Creep and Stress-Relaxation Testing” in this Volume. For metals and

ceramics, creep occurs in the temperature range of about 0.3 T

to 0.6 T

. For polymers, creep deformation is a

factor at temperatures above the glass transition temperature, T

, of a polymer.

Hot Tension and Compression Testing

Dan Zhao, Johnson Controls, Inc.; Steve Lampman, ASM International

Effects of Temperature

In general terms, the effects of temperature on the mechanical behavior of metals can be classified into three

basic ranges based on the application temperature, T, relative to the melting point, T

, of a metal as follows:

• Cold working applications, T < 0.3 T

• Warm working applications, 0.3 T

< T < 0.6 T

• Hot working applications, T > 0.6 T

These general temperature ranges are based on the underlying physical processes that influence mechanical

behavior at different temperatures. For example, the warm working temperature range (0.3 T

< T < 0.6 T

) is

the region of creep deformation. It is also the region of recovery and recrystallization. This includes high-

temperature applications of structural materials (such as those listed in Table 1) and room-temperature testing

of metals with low melting points.

Table 1 Typical elevated temperatures in engineering applications

Application Typical

materials

Typical

temperatures

Homologous

temperatures

T/T

Rotors and piping for steam turbines Cr-Mo-V steels 825–975 0.45–0.50

Pressure vessels and piping in

nuclear reactors

316 stainless steel 650–750 0.35–0.40

Reactor skirts in nuclear reactors 316 stainless steel 850–950 0.45–0.55

Gas turbine blades Nickel-base superalloys 775–925 0.45–0.60

Burner cans for gas turbine engines Oxide dispersion-strengthened nickel-

base alloys

1350–1400 0.55–0.65

Source: Ref 1

In general, strength is reduced at high temperatures, and materials become softer and more ductile as

temperature increases. However, the rate and direction of property changes can vary widely for the yield

strength and elongation of various alloys as function of temperature, as shown in Fig. 1. These changes are due

to various metallurgical factors. For example, there is a significant drop in the ductility of 304 stainless steel in

the temperature range of 425 to 870 °C (800–1000 °F). This ductility drop is from the embrittling effect of

carbide precipitation in the grain boundaries.

Fig. 1 Effect of temperature on strength and ductility of various materials. (a) 0.2 offset

yield strength. (b) Tensile elongation. Source: Ref 2

Other factors, which often cannot be easily predicted, can also affect mechanical behavior at high temperatures.

For example, resolutioning, precipitation, and aging (diffusion-controlled particle growth) can occur in two-

phase alloys, both during heating prior to testing and during testing itself. These processes can produce a wide

variety of responses in mechanical behavior depending on the material. For example, Fig. 2 shows the effect of

exposure time on the high-temperature yield strength and elongation of a precipitation-hardening aluminum

alloy.

Fig. 2 Effect of exposure time on (a) yield strength and (b) elongation at testing

temperature for an aluminum alloy 2024. Source: Ref 2

Effect of Temperature on Deformation and Strain Hardening. As temperature increases, the strength of a

material usually decreases and the ductility increases. The general reduction in strength and increase in ductility

of metals at high temperatures can be related to the effect of temperature on deformation of the material. At

room temperature, plastic deformation occurs when dislocations in the material slip. The dislocations also

intersect and build up in the material as they slip. This build-up of dislocations restricts the slip, and, thus,

increases the forces necessary for continued deformation. This process is known as strain hardening or work

hardening.

At elevated temperatures, dislocation climb comes into play as another deformation mechanism. Further, the

build-up of strain energy from strain hardening can be relieved at high temperatures when crystal imperfections

are rearranged or eliminated into new configurations. This process is known as recovery. A much more rapid

restoration process is recrystallization, in which new, dislocation-free crystals nucleate and grow at the expense

of original grains. The restoration processes can be greatly enhanced by the increase in the thermal activity and

mobility of atoms at higher temperatures. As a result, lower stress is required for deformation, as shown in the

stress-strain diagrams of several materials at elevated temperatures (Fig. 3 4 5 6 7 ).

Fig. 3 Elevated-temperature stress-strain curves in tension for Fe-18Cr-8Ni (Type 301)

stainless steel. (a) 0.508 mm (0.020 in.) sheet full hard from 40% reduction (data average

of longitudinal and transverse). (b) 0.813 mm (0.032 in.) sheet full hard with stress relief

at 425 °C (800 °F) for 8 h. Source: Ref 3

Fig. 4 Effect of high (1.0 s

-1

) and low (0.05 × 10

-4

-1

) strain rates and temperature on

stress-strain curves of 1020 hot-rolled carbon steel sheet (1.644 mm, or 0.064 in.). Source:

Ref 4